ML19317F334
| ML19317F334 | |
| Person / Time | |
|---|---|
| Site: | Oconee  |
| Issue date: | 04/18/1973 |
| From: | Stella V US ATOMIC ENERGY COMMISSION (AEC) |
| To: | Deyoung R US ATOMIC ENERGY COMMISSION (AEC) |
| References | |
| NUDOCS 8001100709 | |
| Download: ML19317F334 (9) | |
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(JR 16 G73 R. C. DeYoung, Assistant Director for Pressurized Water Reactors, L CLAD COLLAPSE TIME CALCULATION, OCONEE 1, DOCKET No. 50-269 In our evaluation of the effects of fuel densification for the Oconce 1 operation, we have reviewed the B&W method to calculate the expected fuel rod collapse time as described in B&W topical reports BAW-10054 and BAW-1387. In addition, we discussed the subject with B&W and Duke Power Company during meetings in the past two months.
Based on our review, we have concluded that the B&W collapse time model is not acceptable. A decailed description of the technical bases for our con-clusion is enclosed and summarized below:
1.
The in-pile test data, on which the semi-empirical B&W method is based, were limited to 3000 hours0.0347 days <br />0.833 hours <br />0.00496 weeks <br />0.00114 months <br /> of exposure, and a significant extrapolation is required for reactor use.
2.
Most of the test rods did not collapse and the collapse time was estimated usinb the consputer code CRECOL, using aa out-of-pile creep law and a critical ovality concept.
3.
The Larson-Miller Parameter was used to determine creep collapse times beyond the range of test parameter data, thus requiring a second extrapolation.
Using the B&W collapse time model together with the conservative assu=ption of no ir. crease of rod internal pressure due to fission gas release, an expected collapse time of 23,400 hours0.00463 days <br />0.111 hours <br />6.613757e-4 weeks <br />1.522e-4 months <br /> is computed. We have used the staff's
- evaluation model, BUCKLE, with the same design and operating parameters as B&W, and; calculated a collapse time of approximately 20,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />.
E c appifcant has stated in BAW-10054 and BAW-1337 that the E&W collapse time evaluation needs to be considered for the first cycle of operation of Oconee 1 only, which is approximately 7,500 effective full power hours.
B&W recently indicated, however, that the first cycle for the upcoming B&W plants (Three Mile Island 1, Oconee 2 and 3) using essentially the same fuel as Oconee 1 will be approximately 12,000 effective full power hours and that they do not expect any' collapse during that time based on their calculation.
We also conclude that no collapse will take place for the Oconee 1 type fuel under Oconee 1 type operating conditions for approximately 12,000 effective full power hours, however, we do not reach that conclusion on the basis of E
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R. C. DeYoung the adequacy of the B&W collapse time evaluation model. Our conclusion is based on the conservative BUCKLE evaluation model and on the experience that was gained at Point Beach 1.
This plant has operated without observed clad collapse of pressurized fuel pins of similar design and under similar conditions for 13,000 effective full power hours.
Therefore, we vill request that B&W revise their present collapse time model by using experi-inental data from longer exposure times and by developing an acceptable ana-lytical approach. This revision, however, should not hold up our present review of fuel densification effects for Oconee 1.
We have informed B&W and Duke Power Company of our position regarding collapse time during our last meeting on April 6,1973, and told them that we intend to reflect this position in our Safety Evaluation Report on fuel densifica-tion. In order to avoid possible delays in the upcoming reviews for other B&W plants, they requested that we provide them with this position in writing r.nd we believe that B&W should be so informed by letter from RP.
Drf-inn! S!;:ned ic p'htor S:ello Victor Stello, Jr., Assistant Director for Reactor Safety Directorate of Licensing
Enclosure:
As Stated cc w/ enc 1:
J. Hendrie, L:TR F. Schroeder, L:TR A. Schwencer, L:PWR-4 T. Novak, L:RSYS D. Ross, L:CPB D. Davis, L:PWR H. Schierling, L:PWR R. Lobel, L:CPB S. Kim, L
- 1. Peltier, L L:PWR-2[.
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An Evaluttion of thn Babcock & Wilcox's Analytical Model to Predict Cladding Collaspe Time 1.
Me thodology :
The creep collaspe time may be predicted by either an empirical or an analytical method. The empirical method is defined here as some kind of curve or expression of the creep cellaspe time, T er' as a functior. of wall thickness, radius of tub 2, cladding tempera-ture, dif ferential pressure and f ast flux as shown below:
T
= F (t, r, T, Ap, 4 )
cr F
An analytical method uses a functional expression of the cladding behavior as given below:
G (w, r, t ;T,$,w, T) = AP F
o t
Shere w denotes the radial displacement of the tube wall and w is o
the given initial ovality.
Some analysts do not use ovality as a parameter in the equation as will be explained later.
Ove rwhelming preference is placed on the analytical method rather than the empirical method because it has the following advantages:
a.
It requires less data points to verify the model.
b.
Extrapolation is more reliable, c.
Usually the equation is given in the form of a differential equation and is less sensitive to input uncertainty such as pressure variation.
Disadvantages for the analytical model are:
a.
It is difficult to construct and requires advanced technique.
b.
It of ten requires assumptions for some basic items such as constitutive equations.
The B&W method is based on a combination of analytical and empirical methods.
They performed a collaspe test in pile with B&W standard Zircaloy tubing. Most of the rods did not collaspe, and they only measured ovality change as a function of time. The maximum time change was less than 3,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />. Because they did not observe the collaspe time, a computer code, CRECOL, was used to predict what they call critical ovality, t.
ovality which corresponds to collaspe time. The v
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use of critical ovality is discussed in Section.2.
Since the test was limited to 3,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />, an extrapolation was performed to go beyond the time range.
In other words, they used a somewhat higher stress. level in the cladding-to shorten the test time and then extrapolated to the lower stress ranges in the teactor by means of the Larson and Miller Parameter. This is another item', in' addition to the use of critical ovality, that we do not feel is adequately
- substantiated.
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There have been two independent approaches taken to arrive at a critical buckling time. One depends on a geometric imperfection
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to provide a self aggravating bending stress which in time produces In this category, we can cite unrealistically large deformations.
The second method Hof f [1], Wilson [2], Merckx [3], and Pan (4].
The is: based on a stability analysis of a differentical equation.
structural material behavior depends on a creep equation and in this case, no initial imperfection is required. The mechanical properties change with time until a small external perturbation results in un-Several authors such as Yamamoto [5], Griffin bounded deflections.
A clear distinc-
[6], and Gerard [7] followed the second approach.
tion between the two methods was made by Jahsman '[8].
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The analytical models of all applicants, including the staff's BUCKLE code so f ar have followed the first method.
Some of the reasons are that the second method is somewhat alien to most engineers and the required material properties or isochronous There are two major difficulties stress-strain curves are scarce.
associated with the first method, however. ' As the initial the imperfections (ovality) decreases and approaches zero in the limit, calculation be'comes singular as shown in Figure 1.
It is difficult to tell at which point the solution becomes unrealistic. The other diffi-culty is that, in certain cases, the ovality change fails to show a clear trend for the instability as time increases. This is indicated in Figure 2.
Therefore, one is faced with an arbitrary choice as to what we mean by a critical time (as collaspe time). Wilson [2] and
!!crckx [3] seems to have this difficulty. Hoff [1] and Pan (4) are the exception as they show a clear trend for the critical time as shown in Figure 3.
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What all this means is that the calculation of ovality with respect to time is quite arbitrary. This is not surprising, since in the eigenvalue problem the deflection has always been arbitrary. That is, one can cal-culate a buckling load without calculating deflection explicitly.
Also note that the ovality calculation was used only indirectly in the time dependent calculations since in the true sense collaspe time corresponds to an infinite ovality or close to it.
In other words,
the ovality value should never be used explicitly in determining collapse time but rather its rate of change, i.e., as it approaches infinity. B&W used the ovality-critical time curve to construct the pressure-Larson-Miller Parameter (LMP) curve.
Since these ovalities are quite arbitrary one can select another set of ovalities and con-struct somewhat dif ferent curves with respect to the LMP whose end consequences are quite large.
The B&W ovality calculation was base on CRECOL, a code developed by d
Butterman, Gelman and Hopkins [10].
By talking to one of the authors, Hopkins, and also reading the reference it was not clear what is meant by critical ovality.
Hence, it was concluded that the use of critical ovality is not recommended.
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The Larson-Miller Parameter (LMP)
The.Ile parameter was originally developed for creep and rupture tests.
The creep time to a certain strain or rupture time depends on stress, temperature and material. To make the interrelationship more manage-able, one can devise a parameter which is some kind of combination of temperature,, time and material such that a single curve can be con-structed, namely rupture time vs. the parameter. This eliminates the need for many curves of stress vs. rupture time at different tempera-tures.
In order for this method to have a meaning the parameter has to'be constant foi any combination of those three independent variables.
Larson and Miller [11] proposed the following relationship:
LMP = (c+ log t)T Where c, t and T are the material constant, time and temperature respectively.
It was found that c is fairly constant for different 1.
materials. Once this was established, then an extrapolation is possible outside of the test range by:
LMP = (c+ log t ) T1 = (c+ log t )T2 t
2 This has great advantages since one can test at a higher temperature Ti which results in a shorter time ti and then extrapolate to t2 for n' lower temparature environment T. The extrapolation is performed of 2
course under a constant stress.
A survey 'f creep by Robatnov [12) states that the range of interpolation i
.iecd limited. He further notes that "In the present state of knovL q.;a, creep calculations on products can only be considered reliable provided that direct experi-mental creep-results are available for the material under consideration throughout the working range of stress and temperature.
Indeed, many investigators are :using the parameter not for extrapolation, but inter-polation over quite a narrow temperature range."
The above discussion has been limited to the use of the parameter on creep and rupture. The use by B&W of the parameter for the collaspe time itself is something entirely new, and its validity can not be establisbod without direct support from actual data and further research.
B&W not onlyj eses the parameter in relation with such a questior.able quantity as critical ovality, they further e.xtrapolate the parameter itself 5eyond the experimental stress range.
In other words, it is an extrapolation on an extrapolation, neither of which.was substcntiated in any way'.
B&W refers to a reference, a report by Woods [13), without further elaboration on the report.
I talked with Mr. Bingman who is one of the co-authors of the report, on the telephone. He said that the t
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parameter was used only for convenience (note that it eliminates the need for many graphs), and he had no intention of extrapolating it.
He also added that he is not aware of anyone usinr, the parameter to predict collaspe time.
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.m Re f e rences :
[1] Ilof f, 'N.T., Jahsman, W.F., and Nadibar, W., "A Study of Creep Collaspe of a Long Circular Cylindrical Shell Under Uniform External Pressure" Journal of the Aerospace Science, Vol. 26, No. 10, pp. 663-674.
[2] Wilson, W.K., "A Method of Analysis for the Creep-Buckling of Tubes Under External Pressure" WAPD-TM-956, Oct. 1970.
[3] Merckx,K. R.2. " Cladding Collapse Calculational Procedures" JN-72-23, Jersey Nuclear Co. Publication (proprietary).
[4] Pan, Y.S., " Creep Buckling of Thin Walled Circular Cylindrical Shells Subject to Radial Pressure and Thermal Gradients" ASME Publication Paper No.170-WA/APM-8.
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[5] Yamamoto, Y., " Considerations of a Stability Criterion for Creep Buckling" J. Mech. Phys. Solids 18, 165-177 (1970).
[6] Grif fin, D.S., " Deformation and Collaspe of Fuel Rod Cladding Due to External Pressure" WAPD-TM-591 Jan. 1967.
[7] Cerard, G., " Theory of Creep Buckling of Perfect Plates and Shells"
-J. of Aero Space Science, Sept. 1962.
[8] Jahsman, W.E., " Creep Stability of a Column with Coupled Geometric Imperfection and Material Behavior Effects" Symposium on Creep in Structure, Springer,1972.
-[9] Pankaskie, P.J., " BUCKLE - An Analytical Computer Code for Calculating Creep Buckling of an Initially Oval Tube".
BNWL-B-253, March 1973.
Buttemer, der., Gelman, A., Hopkins, H.C:., "CRECOL - A Computer
[10)
Program for Estimating Creep Deformation of Slightly Oval Tubes" Gulf General Atomic GAMD-9623, Aug.1969.
[11) Larson, F.R., Miller, J., "A Time - Temperature Relationship for Rupture and Creep Stress" Trvas. ASME, 74 (1952), P765-775.
[12] Robatvov, Yu.N., " Creep Problems in Structural Members" American Elsevier, New York, 1969.
[13) Berman, R.M., Bingman, F.O., et al and edited by Wood, " Properties of Zircaloy-4 lubinS" WAPD-TM-585, Dec.1966.
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